[Retracted] Performance Analysis of High-Speed Deep/Shallow Recessed Hybrid Bearing

Abstract

The present paper proposes a theoretical analysis of the performance of deep/shallow recessed hybrid bearing. It is intended that, on the basis of the numerical results drawn from this study, appropriate shallow recess depth and width can be determined for use in the bearing design process. By adopting bulk flow theory, the turbulent Reynolds equation and energy equation are modified and solved numerically including concentrated inertia effects at the recess edge with different depth and width of shallow recess. The results indicate that the load capacity, drag torque increases as the depth of shallow recess is shallower and the width ratio (half angle of deep recess versus half angle of shallow recess) is smaller. In contrast, the flow rate decreases as the depth of shallow recess is shallower and the width ratio is smaller. Nevertheless, the appropriate design of the depth and width of shallow recess might well induce the performance of high-speed deep/shallow recessed hybrid bearing.

1. Introduction

The hybrid journal bearings provide superior performance characteristics such as minimum friction, large fluid film stiffness, good damping characteristics, and smooth relative motion. Due to the characteristics of excellent stiffness and damping, good rotational accuracy, low friction, and long life, they can be an attractive alternative to rolling bearings in high-speed machine tool spindle [1, 2]. The hybrid bearings have many types, the deep/shallow recess bearings are a common form of hybrid bearings, and the hydrodynamic effects of these bearings are greatly improved and have the merits of larger load capacity, higher stiffness, and superior stability than many other types of hybrid bearings. Despite all that, the main drawback of these bearings is large heat generation in high-speed condition; therefore, the low-viscosity fluid is adopted in the high-speed deep/shallow recess hybrid bearing to satisfy the demands of high-speed machine tool spindle since the low viscosity can significantly reduce the friction loss and heat generation.

During the last few decades, the theoretical analysis of the hybrid bearings with deep/shallow recesses has been carried out and reported in the literature. Among the first studies were the principles of these bearings [3]. Also the static and dynamic characteristics of deep/shallow recessed hybrid bearings were studied, respectively, with different structure of bearing [4–6], different restrictors [7], various recess shapes [8, 9], and considering different flow regime and operational condition [10]. The depth and width of shallow recess were a key design parameter for hybrid bearings with deep/shallow recess. The following review details some important investigation carried out for the influence of shallow recess depth and width on the hybrid bearings.

Chen et al. [11, 12] studied the influence of the shallow recess depth on the performance of hybrid bearing with orifice compensated and capillary compensated. Compared to the conventional deep recess bearing, they put forward the design criterion of shallow recess depth about the hybrid bearing. Helene et al. [13] presented the theoretical analysis of the pressure pattern in a two-dimensional recess of a hybrid journal bearing. They introduced the ratio of the recess depth versus the film thickness. Six values of this ratio are considered, ranging from 4 (shallow recess) to 152 (deep recess), and presented in a systematic manner giving a clear insight of the flow effects intervening in the recess and of their mutual interaction with laminar and turbulent conditions.

The above references have been instrumental to the present development. Due to the high surface speed, the high-speed hybrid bearing with deep/shallow recesses using the low-viscosity fluid is always operated in the turbulent flow regime; meanwhile the inertia effects are also pronounced, especially the concentrated inertia effects at the edge of deep recess. To the best knowledge of authors, a thorough scan of the available literature reveals that there does not exist any comprehensive study related to the performance of the high-speed deep/shallow recessed hybrid bearing with different depth and width of shallow recess considering the turbulent flow regime and inertia effects in thermohydrodynamic lubrication analysis.

In the present study, an analysis is conducted and solutions are provided for the effects of the depth and width of shallow recess on the performance of the high-speed deep/shallow recessed hybrid bearing. By adopting bulk flow theory, the turbulent Reynolds equation and energy equation are modified and solved numerically including concentrated inertia effects at the deep recess edge with different eccentricity ratios.

2. Theoretical Analysis

2.1. The Geometry of High-Speed Hybrid Bearing with Deep/Shallow Recesses

Figure 1 shows the geometry of high-speed hybrid bearing with four deep/shallow recesses and feed orifices. The deep/shallow recess is comprised of two parts, the shallow recess and deep recess; the pressurized fluid is supplied into the deep recess through orifice, a flow from deep recess into the shallow recess and film land.

Figure 1 

Geometry of the high-speed hybrid bearing with deep/shallow recesses.

2.2. Turbulent Reynolds Equation

The generalized nondimensional Reynolds equation for the turbulent flow regime can be written as 𝜕𝐺𝜕𝜃^𝑥𝐻³𝜇𝜕𝑝+𝐷𝜕𝜃𝐿²𝜕𝜕𝑧𝐺_𝑧𝐻³𝜇𝜕𝑝𝜕𝑧𝜕=Λ𝐻𝜕𝜕𝜃+𝜎𝐻,𝜕𝜏 (1) where nondimensional parameters are defined by 𝑥𝜃=𝑅,𝑧𝑧=,𝐿/2ℎ𝐻=𝑐,𝑝𝑝=𝑃_𝑠,𝜇𝜇=𝜇₀,𝜏=𝜔_𝑝𝜇⋅𝑡,Λ=⁰⋅𝜔⋅𝑅²2𝑝_𝑠𝑐²,𝜇𝜎=⁰⋅𝜔_𝑝⋅𝑅²𝑝_𝑠⋅𝑐², (2) where the turbulent coefficients 𝐺_𝑥 and 𝐺_𝑧 depend on the fluid velocity field and are smaller than 1/12 which is the value they take in the case of laminar flow.

2.3. The Bulk Flow Model for Turbulent Lubrication

Returning to the Navier-Stokes equations for a turbulent thin film and disregarding the inertia terms, after integration, the bulk-flow speed components can be expressed in the following form [14]: 𝑉_𝑥=−𝐺_𝑥ℎ²𝜇𝜕𝑝+𝑈𝜕𝑥2,𝑉_𝑧=−𝐺_𝑧ℎ²𝜇𝜕𝑝.𝜕𝑧 (3)

Here 𝑉_𝑥 and 𝑉_𝑧 are the mean values across the film of the stochastic averages of the fluid velocity components. Following the bulk flow theory of Hirs, the pressure gradient depends on the bulk flow velocity [15]: 𝜕−ℎ𝑝=𝜇𝜕𝑥ℎ12𝑘_𝐽+𝑘_𝐵𝑉_𝑥+𝑘_𝐽𝑈2,𝜕−ℎ𝑝=𝜇𝜕𝑧ℎ12𝑘_𝐽+𝑘_𝐵𝑉_𝑧, (4) where the turbulent shear parameters at shaft and bearing (𝑘_𝐽, 𝑘_𝐵) are defined in terms of the flow Reynolds numbers relative to the rotating shaft and stationary bearing surfaces: 𝑘_𝐽=0.066Re_𝐽^0.75,𝑘_𝐵=0.066Re_𝐵^0.75. (5)

The flow Reynolds numbers relative to the rotating shaft and stationary bearing surfaces are Re_𝐵=𝜌ℎ𝜇𝑉²_𝑥+𝑉²_𝑧^1/2,Re_𝐽=𝜌ℎ𝜇𝑉_𝑥−𝑈²+𝑉²_𝑧^1/2. (6)

Comparing (3), (4) the turbulent coefficients are obtained as 𝐺_𝑥1=min;122𝑉_𝑥−𝑈𝑘_𝐽+𝑘_𝐵𝑉_𝑥−𝑘_𝐽𝑈,𝐺_𝑧1=min;212𝑘_𝐽+𝑘_𝐵. (7)

2.4. The Concentrated Inertia Effects at the Edge of Deep Recess

The local acceleration of fluid film from a deep recess into shallow recess and film land causes a sudden pressure drop in the high-speed condition [16, 17]; the pressure at the entrance to shallow recess and film land is modeled from Bernoulli’s equation as 𝑝_𝑒=𝑝_𝑟−𝜌𝑉²_𝑥,𝑧2(1+𝜉)𝑉_𝑥,𝑧⋅𝑛>0. (8)

Note that the entrance loss coefficient 𝜉 is expressed as the empirical (edge) entrance loss coefficients.

2.5. Energy Equation

The energy equation under adiabatic conditions in turbulent flow regime is obtained as follows: 𝜌𝜕𝑐_𝑣𝑇ℎ𝑉_𝑥+𝜕𝑐𝜕𝑥^𝑣𝑇ℎ𝑉_𝑧𝜕𝑧=𝜏_𝑥𝑦0𝑈−𝜏_𝑥𝑦^ℎ₀𝑉_𝑥−𝜏_𝑧𝑦^ℎ₀𝑉_𝑧, (9) where 𝜏_𝑥𝑦0=ℎ2𝜕𝑝+𝜇𝜕𝑥𝑘4ℎ^𝑏𝑉_𝑥−𝑉_𝑥−𝑈𝑘_𝐽,𝜏_𝑥𝑦^ℎ₀𝜇=−ℎ𝑘_𝑥𝑉_𝑥−𝑘_𝐽𝑈2,𝜏_𝑧𝑦^ℎ₀𝜇=−ℎ𝑘_𝑧𝑉_𝑧. (10)

The energy equation becomes in nondimensional form: 𝜌𝜕𝑐_𝑣𝑇𝐻𝑉_𝑥+𝐷𝜕𝜃𝐿𝜕𝑐_𝑣𝑇𝐻𝑉_𝑧𝜕(𝑍)=𝑝_𝑠2𝑇_𝑠𝐻𝑝𝜇𝜕𝜃𝑈+⁰𝑅𝑐²𝑇_𝑠𝜇4𝐻𝑘_𝑏𝑉_𝑥−𝑉_𝑥−𝑈𝑘_𝐽𝑈+𝜇₀𝑅𝑐²𝑇_𝑠𝜇𝐻𝑘_𝑥𝑉_𝑥−𝑘_𝐽𝑈2𝑉_𝑥+𝜇₀𝑅𝑐²𝑇_𝑠𝜇𝐻𝑘_𝑧𝑉_𝑧𝑉_𝑧, (11) where 𝑇=𝑇/𝑇_𝑠.

The fluid viscosity is considered to vary with temperature according to the following formulae: 𝜇=𝑏𝑒^{−𝑎(𝑡−𝑡0)}. (12)

2.6. Drag Torque

The turbulent wall shear stress at the journal surface based on bulk flow theory is given by [18] 𝜏_𝑥𝑦0=ℎ2𝜕𝑝+𝜇𝜕𝑥𝑉4ℎ^𝑥𝑘_𝑏−𝑉_𝑥−𝜔𝑅⋅𝑘_𝐽. (13) The drag torque is given by integration of the wall shear stress at the journal surface: 𝑇_𝑓=^−𝐿/2_−𝐿/20^2𝜋𝑅𝜏_𝑦𝑥𝑅𝑑𝑥𝑑𝑧. (14)

2.7. Mass Conservation at a Deep Recess

The continuity equation at the deep recess is defined by the global balance between the flow through the orifice restrictor (𝑄_in), the recess outflow into the shallow recess and film land (𝑄_out), and the time rate of mass changing due to squeeze effect (𝑄_𝑗). The recess flow continuity equation is expressed as 𝑄_in=𝑄_out+𝑄_𝑗, (15) where the flow through the orifice restrictor is 𝑄_in=𝐴₀𝐶_𝑑2𝑝_𝑠−𝑝_𝑟𝜌. (16)

The recess outflow into the shallow recess and film land is comprised of two parts, the flow across axial and circumferential boundaries of the deep recess: 𝑄_out=𝑄_out𝑎+𝑄_out𝑐. (17)

The time rate of mass changing due to squeeze effect is determined as follows: 𝑄_𝑗=𝐴_𝑟𝜕ℎ.𝜕𝑡 (18)

2.8. Global Energy Balance Equation at a Deep Recess

The energy leaving the boundary of the deep recess must be equal to the energy flowing into the deep recess (Figure 2); that is, 𝐸_in=𝐸_out. (19)(1)The energy flowing into the deep recess is comprised of two parts, the carry-over of hot fluid from upstream to downstream of the deep recess and the fresh fluid from the supply source into the deep recess volume [19]: 𝐸_in=𝑞_in𝑇_𝑠+𝑞_left𝑇_left. (20)(2)The energy leaving the boundary of the deep recess is expressed as 𝐸_out=𝑞_right𝑇_right+𝑞_up𝑇_up+𝑞_down𝑇_down. (21)

Figure 2 

Boundary conditions and conceptual description of global flow rate and energy balance at a deep recess.

2.9. Boundary Conditions

The boundary conditions for the pressure and temperature fields on the high-speed deep/shallow recessed hybrid bearing which are shown in Figure 2 are expressed as follows.

2.9.1. Pressure Boundary Conditions

(1)The pressure at the ends of the bearing is taken to be equal to the ambient pressure. Thus, 𝑝_Γ=Γ2=𝑝_𝑎. (22)(2)The depth of deep recess is much greater than the film thickness; it is assumed that the pressure in each deep recess is uniform and equal to 𝑝_Γ=Γ3=𝑝_𝑟. (23)(3)Fluid inertia at the edges of deep recess is treated through a Bernoulli-type relationship.(4)The supply pressure is equal to 𝑝_Γ=Γ0=𝑝_𝑠. (24)(5)On the 360° extended film land, the pressure field is continuous and single-valued in the circumferential direction: 𝑝(𝜃,𝑧)=𝑃(𝜃+2𝜋,𝑧). (25)

2.9.2. Temperature Boundary Conditions

(1)The temperature of the deep recess and edges is obtained as described above; that is, 𝑇|_ΓΓ=3=𝑇_𝑟,𝑇|_ΓΓ=1=𝑇_𝑟. (26)(2)The supply temperature is equal to 𝑇|_ΓΓ=0=𝑇_𝑠. (27)(3)On the 360° extended film land, temperature field is continuous and single-valued in the circumferential direction: 𝑇(𝜃,𝑧)=𝑇(𝜃+2𝜋,𝑧). (28)

3. Result and Discussion

The parameters of high-speed deep/shallow recessed hybrid bearing are chosen from Table 1.

Table 1 

The parameters of high-speed deep/shallow recessed hybrid bearing.

3.1. Fluid-Film Pressure Distribution

The distribution of fluid-film pressure along circumferential direction at axial midplane is shown in Figure 3. It is observed from the figure that the hydrodynamic effects of high-speed hybrid bearing with deep/shallow recesses are significant and the peak pressure is higher as the width ratio decreases; the shallow recessed bearing (𝑘=0) has the maximum peak pressure; however, the pressure distribution changes slowly and has no peak pressure along circumferential direction for the deep recessed bearing (𝑘=∞) since the hydrodynamic effects are the lowest; meanwhile the pressure has a sharp decrease at the downstream edge of deep recess due to the concentrated inertia effects. Furthermore, from Figure 7(b), it can be seen that the hydrodynamic effects of high-speed deep/shallow recessed hybrid bearing are more pronounced as the depth of shallow recess is shallower. The peak pressure is located at the downstream of shallow recess.

(a)

(b)

Figure 3 

Pressure distribution for the high-speed deep/shallow recessed hybrid bearing.

3.2. Fluid-Film Temperature Distribution

The profile of fluid-film temperature along circumferential direction at axial midplane is depicted in Figure 4. It is noticed that the bearing temperature rise increases as the depth of shallow recess is shallower and the width ratio is smaller due to the increase of drag torque and the reduction of flow rate caused by hydrodynamic effects. The temperature is found higher at almost all the points along circumferential directions for the shallow recessed bearing and has the maximum peak value (almost 60°); however, due to the large flow rate the temperature rise of deep recessed bearing changes not obviously and the value of temperature is not beyond 30° which showed the advantages of hybrid bearing using the low viscosity fluid under the high-speed condition. Further, it is also observed that the temperature rise of high-speed hybrid bearing with deep/shallow recesses is high (almost 60°) with the shallower depth of shallow recess (ℎ_𝑟=10𝜇m) and the smaller width ratio (𝑘=1:4) due to the combination of hydrodynamic and turbulent effects.

(a)

(b)

Figure 4 

Temperature distribution for the high-speed deep/shallow recessed hybrid bearing.

3.3. Static Load Capacity

Figure 5 shows the static load capacity of high-speed deep/shallow recessed hybrid bearing versus the eccentricity ratio. Note that the static load capacity exhibits a remarkable enhancement as the eccentricity ratio is larger owing to the increase of hydrodynamic effects. The static load capacity increases as the depth of shallow recess is shallower and the width ratio is smaller; meanwhile, the static load capacity of deep/shallow recessed bearing is larger than the deep recessed bearing and almost twice as large as deep recessed bearing for smaller width ratio (𝑘=1:4) at large eccentricity ratio (𝜀>0.7). It can be explained that the predominance of hydrodynamic effects gets over hydrostatic effects with the increase of eccentricity ratio. Moreover, when the width ratio becomes smaller (𝑘=1:4) or larger (𝑘=4:1), the static load capacity of deep/shallow recessed bearing is close to the shallow recessed bearing or the deep recessed bearing, respectively.

(a)

(b)

Figure 5 

Static load capacity versus eccentricity ratio.

3.4. Attitude Angle

Figure 6 shows the attitude angle of high-speed hybrid bearing with deep/shallow recesses against the eccentricity ratio. It can be seen that the attitude angle decreases with the eccentricity ratio except for the deep recessed bearing. The attitude angle is larger with the shallower depth of shallow recess and the smaller width ratio especially for small eccentricity ratio due to the increase of hydrodynamic effects. Further, the attitude angle of shallow recessed bearing is the largest (80°) at lower eccentricity ratio (𝜀<0.3) and decreases rapidly with the increase of eccentricity ratio; however, the attitude angle of deep recessed bearing is almost constant at lower eccentricity ratio (𝜀<0.3) and decreases slightly at large eccentricity ratio (𝜀>0.7).

(a)

(b)

Figure 6 

Attitude angle versus eccentricity ratio.

(a)

(b)

Figure 7 

Flow rate versus eccentricity ratio.

3.5. Flow Rate

The eccentricity ratio as a function of the flow rate is shown in Figure 7. It can be found that the flow rate has almost the same value as that of the eccentricity ratio. There is a reduction in the flow rate as the width ratio is smaller since the hydrostatic effects decrease and the flow rate of deep recessed bearing is almost twice as large as the shallow recessed bearing. Moreover, with the increase of eccentricity ratio, the flow rate of deep/shallow recessed and shallow recessed bearing has a tiny increase, while the flow rate of deep recessed bearing decreases slowly. Furthermore, as the depth of shallow recess is deeper, the turbulent effects increase at the shallow recess region and the flow rate also increases which is seen in Figure 7(b).

3.6. Drag Torque

The drag torque of high-speed deep/shallow recessed hybrid bearing as a function of the eccentricity ratio is presented in Figure 8. The drag torque has almost the same value as the eccentricity ratio is lower and slightly increases as the eccentricity ratio is larger (𝜀>0.7). Similarly, The drag torque is larger (twice as big as the deep recessed bearing) with the shallower depth of shallow recess and the smaller width ratio due to the increase of wall shear stress. Furthermore, the drag torque of deep recessed bearing is the smallest (2 N/m) owing to the larger deep recess; contrarily the drag torque is the largest (4 N/m) for the shallow recessed bearing.

(a)

(b)

Figure 8 

Drag torque versus eccentricity ratio.

4. Conclusion

An analysis is conducted and solutions are provided for the effects of the depth and width of shallow recess on the performance of the high-speed deep/shallow recessed hybrid bearing. By adopting bulk flow theory, the turbulent Reynolds equation and energy equation are modified and solved numerically including concentrated inertia effects at the deep recess edge with different eccentricity ratios. The following conclusions can be obtained.(1)The depth and width of shallow recess significantly affect the pressure and temperature distribution of the high-speed deep/shallow recessed hybrid bearing. The pressure and temperature rise of the deep/shallow recessed hybrid bearing increases with the shallower recess depth and the smaller width ratio.(2)The load capacity of the high-speed deep/shallow recessed hybrid bearing significantly increases as well as the flow rate reduces and the drag torque increases with the shallower recess depth and the smaller width ratio especially at large eccentricity ratio. In contrast, the flow rate decreases as the depth of shallow recess is shallower and the width ratio is smaller.(3)In summary, from the view for the high-speed deep/shallow recessed hybrid bearing design, when the eccentricity ratio is large (𝜀>0.4), the depth of shallow recess should be designed shallower (ℎ_𝑟<30𝜇m) and the width ratio should also be designed smaller (𝑘<1:1) for the high-speed deep/shallow recessed hybrid bearing; otherwise, the deeper shallow recess and the larger width ratio are required in order to reduce the friction power consumption and simplify the design and manufacturing.

Nomenclature

ℎ:	Film thickness
𝐻:	Nondimensional film thickness
𝜃:	Angular coordinate
𝜔:	Journal rotating speed
𝜌:	Density of lubricant
𝑝𝑒:	Pressure at deep recess edge
𝐺𝑥, 𝐺𝑧:	Turbulent coefficients of circumferential and axial turbulent coefficients
𝑘𝐽, 𝑘𝐵:	Turbulent shear parameters at shaft and bearing
𝑉𝑥, 𝑉𝑧:	Bulk flow circumferential and axial velocities
𝜏𝑥𝑦, 𝜏𝑧𝑦:	Wall shear stresses
𝜏𝑥𝑦|0:	Turbulent wall shear stress at shaft surface
𝑇𝑅:	Drag torque
𝐴0:	Orifice sectional area
𝑑𝑐:	Diameter of orifice
𝐴𝑟:	Recess area
𝐶𝑑:	Empirical orifice discharge coefficients
𝑄in, 𝑄out:	The flow through the orifice restrictor and the deep recess outflow into the film land
𝑄out𝑐, 𝑄out𝑎:	Circumferential and radial mass flow
𝐸in, 𝐸out:	The energy flowing into the deep recess and leaving the boundary of the deep recess
𝑇𝑠:	Water inlet temperature
𝑇𝑟:	Temperature within deep recess and downstream edge
𝑇left:	Temperature in upstream edge
𝜉:	Empirical recess-edge entrance loss coefficients.

Subscripts

𝑠:	Supply condition
𝑟:	Bearing recesses
𝐵, 𝐽:	Bearing and shaft.

Acknowledgments

The authors gratefully acknowledge the support of the National Science Foundation through Grant nos. 50775036, 11172065 and Jiangsu University-Industry Collaboration Project through Grant no. BY2012199.